NASA TECHNICAL NOTE
UO
CO
wo
NASA TN D-7355
MATERIALS TECHNOLOGY PROGRAMS
IN SUPPORT OF A MERCURY
RANKINE SPACE POWER SYSTEM
by Phillip L. Stone
Lewis Research Center
Cleveland, Ohio 4413 3
NATIONAL AERONAUTICS AND SPACE ADMINISTRATION « WASHINGTON, D. C. • SEPTEMBER 1973
1. Report No, 2. Government Accession No.
NASA TN D-7355
3. Recipient's Catalog No.
4. Title and Subtitle
MATERIALS TECHNOLOGY PROGRAMS IN SUPPORT
OF A MERCURY RANKINE SPACE POWER SYSTEM
5. Report Date
September 1973
6. Performing Organization Code
7, Author(s)
Phillip L. Stone
8. Performing Organization Report No.
E-7382
10. Work Unit No.
501-21
9. Performing Organization Name and Address
Lewis Research Center j
National Aeronautics and Space Administration
Cleveland, Ohio 44135
1 1 . Contract or Grant No.
13. Type of Report and Period Covered
Technical Note
12. Sponsoring Agency Name and Address
National Aeronautics and Space Administration
Washington, D. C. 20546
.......
14. Sponsoring Agency Code
15. Supplementary Notes
16. Abstract
This report summarizes a large portion of the materials technology that was generated in support
of the development of a mercury -rankine space power system (SNAP -8) . The primary areas of
investigation reported are (1) the compatibility of various construction materials with the liquid
metals mercury and NaK, (2) the mechanical properties of unalloyed tantalum, and (3) the devel-
opment of refractory metal /austenitic stainless steel tubing and transition joints. The primary
results, conclusions, and state of technology at the completion of this effort for each of these
areas are summarized in this report. Results of possible significance to other applications are
highlighted.
For sale by the National Technical Information Service, Springfield, Virginia 22151
CONTENTS
Page
SUMMARY 1
INTRODUCTION 2
LIQUID METAL CORROSION STUDIES 3
Mercury Corrosion 4
Conventional alloys 4
Refractory metals 6
Cavitation damage . 8
NaK Corrosion 9
Conventional iron- and nickel-base alloys 9
Refractory metals ll
MECHANICAL PROPERTIES OF UNALLOYED TANTALUM . , i2
Tensile Properties 12
Creep Properties . . 14
Low Cycle Fatigue Tests . . . . 15
FABRICATION AND EVALUATION OF BIMETALLIC TUBING AND JOINTS 17
Initial Investigations . . 18
Tantalum/316 Stainless Steel Tubing Development 20
General . 20
Hot coextruded tubing - mandrel 24
Hot coextruded tubing - filled billet . 24
Explosion welded tubing 25
Evaluation of tantalum/316 stainless steel bimetallic tubing 25
Tantalum/316 Stainless Steel Transition Joint Development 28
General . 28
Brazed joints 29
Hot coextruded joints ................................ 31
Evaluation of joints 32
APPLICABILITY OF RESULTS 34
SUMMARY OF RESULTS 34
APPENDIX - PROCEDURES USED TO COEXTRUDE TANTALUM/
316 STAINLESS STEEL BIMETALLIC TUBING 36
REFERENCES. 44
iii
MATERIALS TECHNOLOGY PROGRAMS IN SUPPORT OF A MERCURY
RANKINE SPACE POWER SYSTEM
by Phillip L Stone
Lewis Research Center
SUMMARY
During the development of the SNAP-8 mercury -rankine space power system, a sig-
nificant amount of support materials technology was performed at the NASA -Lewis
Research Center and at other laboratories under NASA contract. This technology was
primarily concerned with (1) the compatibility of some conventional alloys and refractory
metals with two liquid metals, mercury (Hg) and a sodium -potassium eutectic alloy
(NaK) , (2) the determination of the mechanical properties of unalloyed tantalum (Ta) ,
and (3) the fabrication and evaluation of refractory metal/austenitic alloy (primarily
Ta/316 stainless steel (316 SS)) bimetallic tubing and transition joints.
Asa result of this materials support work, the following conclusions were made:
(1) The refractory metals Ta, niobium-1 -percent zirconium (Nb-lZr), and the Ta alloy
T-lll should not be attacked by liquid Hg at temperatures up to 650° C (1200° F) .
(2) The NaK corrosion rates of the major SNAP -8 primary reactor loop materials,
Hastelloy N and 316 SS, appear to be acceptably low «0.004 cm/10 4 hr or < 0.0015
in./10 4 hr) at temperatures up to 700° C (1300° F), providing that the oxygen level in
the NaK is maintained at less than 30 ppm. (3) Tantalum having a uniformly distributed
oxygen concentration of 115 ppm or less should not be attacked by purified NaK at tem-
peratures up to 730° C (1350° F). (4) Recrystallized fine-grained Ta tubing is more
creep resistant under uniaxial loading than large -grained material at 730° C (1350° F).
(5) Recrystallized Ta bar has a low-cycle fatigue life in excess of 1000 cycles at a plas-
tic strain range of 0.02 centimeter per centimeter (in. /in.) at temperatures up to
730° C (1350° F). (6) Welded butt joints, tee joints, and tube-to-header joints for re-
fractory metal/austenitic alloy tubing can be successfully produced. (7) The most suit-
able refractory metal/austenitic alloy bimetallic combination to be used as tubing for Hg
containment at temperatures to 730° C (1350° F) is Ta/316 SS. (8) A successful fabrica-
tion method for producing Ta/316 SS bimetallic transition joints involves brazing and a
tongue -in -groove design.
INTRODUCTION
This report summarizes a large portion of the materials technology that was devel-
oped in support of a previously planned nuclear, mercury- rankine space-power genera-
tion system termed SNAP-8. (A summary of the SNAP- 8 program, since completed, can
be found in ref. 1. ) Most of this materials technology has been previously reported, but
in so many diverse ways and places and under so many different titles, as to make it ex-
tremely difficult to develop total context or to reach overall conclusions. It is, there-
fore, the primary purpose of this report to summarize the major results of these many
efforts and to give more comprehensive overall conclusions so that future investigators
may have readily available information upon which to build. Although many other mate-
rials investigations in support of SNAP- 8 were performed by Aerojet General Corpora-
tion, NASA’s prime contractor for this system (ref. 1), the work described in this report
was performed either at the NASA-Lewis Research Center or under other NASA con-
tracts. More detailed discussion of each of the studies covered can be found in the refer-
ences cited.
The original goal of the SNAP -8 mercury- rankine system was to provide 35 kilo-
watts of electrical power (kW ) for space applications for at least 10 000 hours. This
V
goal gradually evolved to a 90 kW requirement and a 40 000 hour life. In both cases,
V
thermal energy for the system was provided by a uranium -hydride- fueled nuclear reac-
tor, which was cooled by flowing liquid sodium-potassium eutectic alloy (NaK). The NaK
was pumped to a heat exchanger (boiler) where it transmitted its thermal energy to the
mercury (Hg) working fluid. (See fig. 1 for a SNAP -8 power conversion system sche-
Figure 1. - SNAP-8 power conversion system schematic.
CD-11490-17
2
matic. ) The Hg was boiled and superheated, and flowed to a turbine alternator, which
generated the electrical power. The Hg was then condensed and pumped back to the
boiler. Cycle waste energy was removed from the condenser-heat exchanger to a radia-
tor by means of another pumped-NaK loop for rejection to the space environment.
It was originally assumed that this SNAP -8 system could readily build on existing
technology from the similar SNAP-2 system (mercury -rankine, 3 kW g ), and thus no
significant effort would be required in the materials development area. But, shortly
after the SNAP -8 program began, it was apparent that SNAP -2 technology was generally
inadequate and that the materials area would involve a great amount of technology devel-
opment. The most difficult materials problems were in the realm of Hg containment,
particularly in the Hg boiler. The boiler was first constructed of the cobalt -base alloy
L-605. When serious Hg corrosion of this alloy was observed, an interim change to
9-percent-chromium - 1-percent-molybdenum steel was made. Finally, it was deter-
mined that a refractory metal was required to reliably achieve the desired Hg corrosion
resistance and system life. Unalloyed tantalum (Ta) was chosen. It was believed nec-
essary to isolate the Ta from the flowing NaK stream to prevent it from gettering such
ductility impairing elements as carbon and nitrogen. To this end, two boiler designs
were initiated, both of which are described in more detail later in this report. One de-
sign used tantalum /austenitic stainless steel (SS) bimetallic tubing. The other design
involved the use of a double -containment configuration in which several unalloyed Ta
tubes, each surrounded successively by nonflowing NaK and a stainless steel tube, were
further contained in a stainless steel outer boiler shell within which the primary loop
NaK flowed. This desigii employed a Ta/316 SS transition joint at each end of the boiler .
The materials investigations discussed in this report were therefore primarily con-
cerned with three areas (1) the compatibility of some conventional iron, nickel, and
cobalt-base alloys and some refractory metals with Hg and NaK, (2) the determination
of design -allowable mechanical properties of unalloyed Ta, and (3) the fabrication and
evaluation of refractory metal /austenitic alloy bimetallic tubing and transition joints.
LIQUID METAL CORROSION STUDIES
At various times during the conduct of the SNAP-8 program it became necessary to
determine the compatibility of certain candidate SNAP -8 construction materials with Hg
and NaK. Many of these corrosion studies, concerned with conventional iron, nickel,
and cobalt-base alloys and also refractory metals, were conducted either at NASA-Lewis
or at other laboratories working under contract to Lewis. The studies included capsule
tests, loop tests, and cavitation experiments. The highlights and prime conclusions of
these studies are summarized in the following sections.
3
Mercury Corrosion
Conventional alloys . - Initially, a cobalt base alloy, L-605, was selected as the ref-
erence fabrication material for the SNAP-8 system because of its apparent successful
use in the SNAP -2 system. In order to determine the Hg corrosion resistance of L-605
at system temperatures, reflux capsule tests and a pumped -loop test were conducted at
Lewis (refs. 2 and 3). Several materials other than L-605 were also tested as reflux
capsules. They were 9Cr-lMo steel, the cobalt-base alloy H-8187, and the iron-base
alloys AM 350 and AM 355. No nickel-base alloys were tested because of the poor Hg
corrosion results previously obtained at TRW, Inc. , in support of the SNAP-2 program
(refs. 4 and 5).
The reflux capsule tests were conducted over the temperature range 540° to 700° C
(1100° to 1300° F) for times up to 5000 hours. The results of the tests indicated gross
Hg attack of all materials other than the 9Cr-lMo steel. Comparative photomicrographs
of L-605 and 9Cr-lMo steel are shown in figure 2. The L-605 developed a deep porous
layer, and the 9Cr-lMo attack consisted of only a shallow surface effect.
9Cr-lMo Steel
L-605
Figure 2, - Typical appearance of mercury-corroded areas of reflux cap-
sules. Photomicrographs of longitudinal section of capsule wall.
Capsule inner surface at far right of photomicrograph. Unetched;
test temperature, 590° C (1100° FI; test period, 1000 hours (ref. 2 ).
4
Figure 3. - NASA - Haynes 25 mercury corrosion loop (ref. 3).
The L-605 corrosion loop (fig. 3) was operated for 1147 hours at a peak liquid tem-
perature of 580° C (1075° F) and at an average liquid velocity of 240 centimeters per
second (8 ft/sec). The corroded porous layer thickness observed was 0. 020 to 0. 025
centimeter (0. 008 to 0. 010 in. ). This was approximately 10 times that observed in the
reflux capsule tests for approximately the same test time. The increased corrosion was
attributed to the effect of the higher Hg liquid velocity in the loop test.
Based on the results of the capsule and loop tests, it was concluded that L-605 was
not a good Hg containment material for SNAP -8 use. Thus, it was decided to substitute
9Cr-tMo steel as the SNAP -8 boiler reference material.
Following the decision to change to 9Cr-lMo steel as the boiler reference material,
a contract was awarded to TRW, Inc. , for the construction and operation of a Hg forced
circulation loop to further study corrosion mechanisms in 9Cr-lMo steel and also cor-
rosion product separation techniques (ref. 6) . This loop (fig. 4) was operated for 2918
hours at an average boiling temperature of 580° C (1075° F) and at an average liquid
velocity of 0.61 centimeter per second (0.02 ft/sec). The corrosion of the 9Cr-lMo
steel was insignificant. But in retrospect, it was believed that the low liquid velocity in
the loop, compared with the velocities newly being projected for SNAP -8, rendered the
corrosion results to be of little direct applicability to the new system requirements.
However, the corrosion product separators in the vapor portion of the system removed
about 54 percent of the total corrosion products found in the system, and the separator
in the liquid region removed about 25 percent. Based on this success, it was concluded
that separators could also be effective in a larger system should the need arise.
Some Hg corrosion loop testing of 9Cr-lMo steel was also conducted by Aerojet
General Corporation under the scope of the SNAP-8 prime contract (ref. 7). This test-
ing established conclusively that Hg corrosion of 9Cr-lMo steel was velocity dependent.
Also, at the velocities required in the boiler for stable Hg boiling, the corrosion rate
5
Wmimm Superheater heaters
Figure 4 - TRW 9Cr-lMo steel mercury corrosion loop (ref. 6).
was considered to be unacceptable for the required 10 000-hour system life (i. e. , a uni-
form 0.0025 cm/100 hr or 0.001 in. /100 hr).
Although it was not expected to be more Hg corrosion resistant than 9Cr-lMo steel,
a modified 9Cr-lMo steel was also used for SNAP -8 boiler construction. This material
was considerably stronger than the standard alloy as a result of the addition of very
small amounts of niobium, vanadium, boron, nitrogren, and zirconium and the use of
a 1040° C (1900° F) normalize and 730° C (1350° F) temper heat treatment. The im-
proved strength properties are clearly illustrated in figure 5 in which the modified alloy
in the normalized and tempered condition is compared with standard 9Cr-lMo steel,
304 SS, and 316 SS, all in the annealed condition. The modified alloy indicated no ther-
mally induced instabilities even after 18 000 hours of stress-rupture testing at 650° C
(1200° F) .
Refractory metals . - As a result of the 9Cr-lMo steel Hg corrosion testing, this
material was replaced by unalloyed Ta as the SNAP -8 boiler reference material. The
selection of Ta was based primarily on the minimum solubility of Ta in Hg at tempera-
tures well above the 590° C (1100° F) Hg boiling temperature, as determined at the
6
25 . —
700 800 900 1000 1100 1200
Temperature, °F
Figure 5. - Strength properties of modified 9 Cr-l Mo steel and other conven-
tional materials as function of temperature. Note: Allowable stress values
are from the ASME boiler and pressure vessel code.
Brookhaven National Laboratory under an AEC sponsored program (ref. 8) . The
Brookhaven data are shown in figure 6. When Ta was selected as the SNAP-8 boiler
reference material, it was recognized that other refractory alloys such as niobium -
1-percent zirconium (Nb-lZr) and T-lll (nominal composition, Ta-8W-2Hf) offered
higher strengths and probably improved corrosion resistance to NaK. But the main
deterrent to their use was the uncertainty of their resistance to Hg corrosion, particu-
larly in the strained state. Data from reference 8 had indicated possible problems in
that regard. Since it was possible, however, that it might become necessary to change
to a higher strength refractory alloy, a test program was initiated to ascertain whether
Hg corrosion and/or stress corrosion problems actually existed with Nb-lZr and T-lll.
The test program was conducted under NASA sponsorship by the General Electric
Company (ref. 9). Sheet specimens of Ta, Nb-lZr, and T-lll in the as-bent and in
several as-bent -and-annealed conditions were exposed to liquid Hg isothermally in tan-
talum capsules at 650° C (1200° F) for 1000 hours. All specimens were totally unaffected
by this exposure. It was therefore concluded that either Nb-lZr or T-lll could be sub-
stituted for Ta in the SNAP -8 system should the need arise.
Some Hg corrosion testing of Ta was also conducted by Aerojet General Corporation
under the scope of the SNAP -8 prime contract (refs. 1, 10, and 11). Included in the
testing was an 8700-hour test of a full-size SNAP-8 Hg boiler. General Electric Com-
7
780 727 679 636 596 560 527 496
Temperature, °C
Figure 6. - Solubility of elements in mercury
(ref. .8).
pany tested another full-size SNAP -8 boiler for 15 250 hours (ref. 12). In none of these
tests did the Ta show any sign of Hg attack, thus verifying its acceptability for long-
term service.
Cavitation damage . - Based on early 9Cr-lMo steel corrosion loop testing (ref. 7),
it appeared that certain mechanisms present during the Hg boiling process might re-
semble a cavitation situation. As a result, studies were initiated at Lewis (ref. 13) and
at the University of Michigan (ref. 14) to determine simulated cavitation effects of Hg on
various materials. At Lewis, 9Cr-lMo steel was compared with three materials be-
lieved to have good resistance to cavitation damage due to their strength and/or hard-
ness: L-605, Hastelloy X, and Stellite 6B. The test specimens were vibrated in 150° C
(300° F) liquid Hg at 25 000 hertz at a peak -to -peak displacement amplitude of 0.0045
centimeter (0.00175 in.). The 9Cr-lMo steel was the least cavitation resistant of the
materials tested (see fig. 7). In the University of Michigan study, Ta was compared
with 9Cr-lMo steel and niobium in liquid Hg at 260° C (500° F). The frequency of vi-
bration of the test specimens was 20 000 hertz, and the amplitude was 0.05 centimeter
(0.002 in.). It was determined that Ta was considerably less resistant to cavitation
damage than was the 9Cr-lMo steel.
8
Annealed 9Cr-lMo Steel (1 hr)
Hastelloy X (1 hr)
L-605 11 hr) Stellite 6B (2 hr)
Figure 7. - Damaged surfaces of specimens after exposure to cavitation in mercury at 150° C
(300° F) (from ref. 13).
Based on the results of these tests, it was concluded that neither 9Cr-lMo steel nor
Ta would be very resistant to cavitation should such a phenomenon actually be present
in a SNAP -8 boiler. But the subsequent Ta corrosion testing revealed no evidence of
cavitation damage (refs. 1 and 10 to 12) . Thus it was concluded that either no cavitation
situation actually existed in the boiler or that cavitation effects were greatly exaggerated
in the laboratory simulation tests.
NaK Corrosion
Conventional iron- and nickel -base alloys . - While 9Cr-lMo steel was the reference
material for SNAP -8, an engineering evaluation of the compatibility of the primary loop
constructional materials with the NaK reactor coolant was performed . The main goal
was to determine where the maximum corrosive attack would occur and how severe it
would be. Asa result, a NaK corrosion loop program was contracted to the Oak Ridge
National Laboratory (ORNL) (ref. 15). Eleven multimaterial loops, constructed of
9
Figure 8. - Diagram of Oak Ridge National Laboratory NaK corrosion loops <ref. 15).
tubular sections of 316 SS, 9Cr-lMo steel , the nickel-base alloy Hastelloy N, and
347 SS (fig. 8) , were operated for times ranging between 700 and 5200 hours at maxi-
mum NaK temperatures of 700°, 760°, or 790° C (1300°, 1400°, or 1450° F). The
700° C (1300° F) loops were of the greatest interest because they were expected to be
the most representative of the actual conditions at the SNAP -8 reactor outlet. The NaK
used was reactor grade.
Corrosion of Hastelloy N (the material that was expected to be the most adversely
affected in the SNAP- 8 primary loop, based on previous ORNL data from sodium experi-
ments) was less than 0.004 centimeter per 10 4 hours (0.0015 in. /10 4 hr) at 700° C
(1300° F) and at a NaK oxygen level of less than 30 ppm. This was believed to be ac-
ceptably low for a long-life system. At the same NaK oxygen level, the iron-base alloys
exhibited minimal corrosive attack. A NaK oxygen level of 80 ppm did not noticeably
change the corrosion rate of the Hastelloy N. The iron -base alloys, however, were sig-
nificantly adversely affected at this higher oxygen level. One deleterious effect of NaK
exposure was the decarburization of the 9Cr-lMo steel and concurrent carburization of
the 300-series stainless steels. This loss of carbon reduced the 1000-hour stress rup-
ture strength of the 9Cr-lMo steel by about 40 percent.
In addition to the corrosion portion of the program, ORNL was to provide informa-
tion on the behavior and control of the hydrogen (from the reactor fuel) present in the
NaK and on the diffusion of hydrogen from the primary loop into the power conversion
Hg loop. From these hydrogen investigations, it was concluded that NaK hydride would
10
precipitate in the NaK loop at a temperature of 160° C (320° F). It was also determined
that the level of hydrogen in the NaK loop could be significantly reduced by means of a
2. 5-percent bypass flow through a 130° C (260° F) cold trap, or by the use of a small
quantity (about 0. 1 wt. %) of lithium in the NaK to getter the hydrogen. The resultant
hydride could then be effectively cold trapped at a higher temperature (i. e. , 200° to
260° C or 400° to 500° F). The permeability of the primary loop materials to hydrogen
was also experimentally determined.
Refractory metals . - It is known that oxygen-contaminated Ta is subject to corro-
sive attack by NaK at elevated temperatures. The extent and character of the attack is
dependent both on the bulk oxygen level and on the distribution of the oxygen in the Ta.
Since there was a considerable amount of Ta in contact with NaK (albeit nonflowing NaK)
in the double containment boiler, it was important to determine quantitatively the extent
and character of NaK corrosion over the probable range of oxygen levels and oxygen dis-
tributions in the Ta.
This contracted test program was conducted by the General Electric Company
(ref. 16). Tantalum specimens were purposely contaminated to achieve homogeneous
oxygen concentrations of 115, 220, 270, or 520 ppm. Additional contaminated and un-
contaminated specimens were gas tungsten-arc (GTA) welded in pure helium or helium
contaminated with air to evaluate the combined effects of welding, welding gas purity,
and preweld oxygen concentration of the Ta on the corrosion resistance of the Ta to NaK.
The specimens were subsequently exposed to reactor grade NaK at 730° C (1350° F) for
1000 hours in isothermal capsule tests to determine the threshold oxygen concentration
for corrosion. Some specimens were also exposed at 650° C (1200° F) for 100 hours to
determine the effects of temperature and time on corrosion.
The two major conclusions of this test program were as follows: (1) Tantalum hav-
ing a uniformly distributed oxygen concentration of about 115 ppm or less will not be at-
tacked by NaK at temperatures up to 730° C (1350° F), but attack on the Ta will definite-
ly occur at a uniformly distributed oxygen concentration of 270 ppm and above (fig. 9(a)).
(2) Gas tungsten arc welding of contaminated Ta specimens changed the morphology of
the subsequent NaK corrosion; that is, the corrosion generally was worse in the weld
and heat affected zone than in the base metal (fig. 9).
MECHANICAL PROPERTIES OF UNALLOYED TANTALUM
When the SNAP -8 double containment boiler was being designed, very little tensile
data and virtually no long-time creep or low-cycle fatigue data were available for un-
alloyed Ta at 730° C (1350° F) and below. Therefore, a series of mechanical property
tests were conducted to obtain these data. AH tests were performed under contract to
Lewis. The results are summarized in the following sections.
Tensile Properties
The Ta tensile testing program was conducted by Metcut Research Associates.
Specimens from the actual lots of tubing, plate, sheet, and bar that were to be used in
the fabrication of the boilers were tested. Uniaxial tensile specimens were machined
from longitudinal elements of the seamless boiler tubing and were tested with special
grips designed to accommodate the tube curvature. All elevated temperature tests were
-4 fi
conducted in a vacuum of <6.7X10 newton per square meter (5x10 torr) ; most of the
specimens were tested at 730° C (1350° F). Results of the tests are presented in refer-
ences 17 and 18. Table I is a summary of these data.
12
TABLE I. - TENSILE PROPERTIES OF UNALLOYED TANTALUM IN THE
RECRYSTALUZED CONDITION
Tantalum material
Temperature
Tensile strength
Yield strength
Elongation,
percent
Reduction
of area,
percent
°c
0
F
MN/m 2
ksi
MN/m 2
ksi
0.396 cm (0.156 in.)
730
1350
119
17.3
48
7.0
56
93
sheet
119
17.3
45
6.6
52
95
121
17.5
41
5.9
61
92
121
17.5
45
6.6
56
88
121
17.6
52
7.6
51
95
l
125
18. 1
60
8.7
55
91
0.409 cm (0.161 in.)
9
1350
122
17. 7
68
93
sheet
9
1350
118
17.1
61
89
mm
1350
119
17.3
.
74
86
0.635 cm (0.250 in. )
730
1350
n
15.5
43
6.2
64
90
plate
—
15.9
37
5.3
62
79
1 I
16.0
34
5.0
44
79
BB
16.3
34
5.0
61
85
S|
16.8
53
7.7
52
73
1
17.5
34
5.0
59
84
2. 54 cm (1.00 in.)
m
mm
18.0
74
10.7
32
81
plate
13
BB
21.0
62
9.0
41
84
S9
BB
21.2
58
8.5
37
63
2.54 cm (1.00 in.)
mm
21.7
66
9.6
44
84
plate
1
22.8
67
9.7
32
82
1
EM
BB
23.4
64
9.3
40
79
1.65-cm (0.652-in.)
730
1350
161
23.4
61
8.8
34
a 63
i.d. by 0. 13-cm
136
19.7
64
9.3
45
b 71
(0.051 -in.) wall
157
22.8
73
10.6
38
c 67
176
25.5
70
10.1
44
C 59
199
28.8
88
12.8
33
d 65
179
25.9
48
7.0
38
e 70
139
20.2
58
8.4
47
*74
161
23.4
59
8.6
40
f 67
160
23.2
52
7.6
35
f 67
154
22.4
66
9.6
39
f 68
595
1100
181
26 . 2
68
9.9
27
73
595
1100
179
26.0
56
8.2
25
74
425
200
29.0
87
12.7
37
77
425
9
211
30.6
87
11.8
33
76
260
B
218
31.6
70
10 . 1
45
79
260
mm
212
30.8
68
9.9
45
77
25
75
278
40.3
170
24.6
58
74
25
75
267
38,8
125
18.1
54
80
a Heat A, c Heat C. e Heat E.
b Heat B. d Heafc D. f Heat F.
13
340, —
0 200 400 600 800 1000 1200 1400
Test temperature, °F
Figure 10. - Ultimate tensile and yield strength of unalloyed recrystallized tantalum tubing.
Figure 10 is a plot of the ultimate tensile and yield strength data for the tubing
(averaged). At 730° C (1350° F) the most extensive scatter in the test results for a given
shape was observed for the tubing specimens (see table I). This may have been the re-
sult of the slight chemistry variations from heat to heat and/or the grain size variations
(ASTM grain size No. 3 to 7) from tube to tube. The greatest scatter in data occurred
among the different shapes, that is, tubing, plate, sheet, and bar. This was expected
because the thermomechanical history for each shape was different and this is known to
noticeably affect the mechanical properties of most materials. All materials tested
showed adequate tensile properties for their intended applications in the boiler. The goal
for the tubing was 0. 2 -percent yield strength and an ultimate tensile strength about as
high as 9Cr-lMo steel at 730° C (1350° F), that is, 55 and 110 meganewtons per square
meter (8.0 and 16.0 ksi), respectively.
Creep Properties
A uniaxial creep testing program was conducted by TRW, Inc. (ref. 19). The high-
est stresses on the Ta, as determined by a stress analysis of the boiler, existed in the
Ta tubing and in the Ta dome -shaped manifold. Therefore, most of the creep specimens
were machined from the actual lots of tubing (1. 65 -cm (0. 652-in. ) i. d. and 0. 13 -cm
(0.051 -in. ) wall thickness) to be used in the boilers or from the actual sheets
(0. 41 -cm (0. 16-in.) thick) from which the manifolds were to be formed. A sheet speci-
14
men was also tested in a prestrained condition to simulate the actual manifolds, which
are strained 35 to 45 percent during fabrication. The tubing specimens were machined
so as to concentrate the stress in the gage area. All tests were conducted in a vacuum
of 1.3x10 newton per square meter (1x10 torr), and most were tested at 730 C
(1350° F) .
The results for the tubing specimens fell into two separate ranges, depending on the
grain size of the test material. (Typical results are shown in fig. 11(a) .) The larger
grained specimens (ASTM grain size No. 3 to 4) tended to be weaker than the smaller
grained specimens (ASTM grain size No. 4 to 7) , which is typical of material tested be-
low its equicohesive temperature. Steady-state creep rate for the small-grained speci-
mens is shown in figure 11(b). The aforementioned prestrained sheet specimen was de-
formed 30 percent before testing, which was the maximum uniform elongation that could
be achieved. This was considered to be an adequate first-order approximation of the
condition of the material after forming into the dome-shaped manifold. The results of
this test revealed a drastic difference between the recrystallized and prestrained mate-
rial. The time to 1 percent creep at 44. 8 meganewtons per square meter (6500 psi) and
730° C (1350° F) for the recrystallized material was 45 hours and for the prestrained
material about 37 000 hours (extrapolated from a 4900-hr test). It is possible that in
service some stress relieving might occur, but the actual stress on the manifold was
not actually expected to be as high as the test point stress. It was concluded, therefore,
that the Ta dome-shaped manifold would have adequate creep strength for its application
in the boiler if used in the prestrained condition. It was also concluded that recrystal-
lized fine-grained Ta tubing (ASTM grain size No. 5, or higher) would have adequate
creep strength for its intended application in the boiler (i. e. , <2 percent in 40 000 hr at
730° C (1350° F) for stresses of about 20. 7 MN/m^ (3000 psi)).
Low Cycle Fatigue Tests
Since the SNAP-8 boiler had to be capable of 100 startups and shutdowns for ground
testing, it was necessary to obtain data on the low-cycle fatigue behavior of unalloyed Ta.
This testing program was conducted at the General Electric Research Laboratory
(ref. 20). Mechanical strain was used to simulate the strain effected by the temperature
cycles of startup and shutdown. Recrystallized 1.27 -centimeter (0, 5-in.) diameter bar
specimens were used. Because of the extreme sensitivity of Ta to environmental con-
tamination at high temperature, a special titanium susceptor -heater in combination with
a highly purified flowing argon gas was used during the testing. Tests were conducted in
a tightly sealed chamber over the temperature range 20° to 760° C (70° to 1400° F) with
emphasis on tests at 320°, 590°, and 730° C (600°, 1100°, and 1350° F). The resulting
15
Stress, MN/m2 Percent creep
A STM grain size No. 3-4
ASTM grain size No. 4-5
A STM grain size No. 6-7
(a) Effect of grain size on creep rate of unalloyed tantalum tubing. Temperature, 730° C (1350° F); stress,
44.8 meganewtons per square meter {6500 psi).
Heat
ASTM grain
size No.
Temperature,
°C <°F)
o
60249
5
730 (1350)
A
60065
6 to 7
730 (1350)
O
60381
4 to 5
730 (1350)
□
60249
5
790 (1450)
Steady-state creep rate, percent/hr
(b) Steady-state creep rate for unalloyed tantalum tubing.
Figure 11. - Creep behavior of unalloyed tantalum tubing.
Cycles to failure
Figure 12, - Low cycle fatigue behavior of unalloyed tantalum bar.
data are shown in figure 12 . Analysis of the data indicated that the low-cycle fatigue life
of unalloyed Ta at the strain ranges to which it would be exposed in the SNAP -8 boiler
(maximum of 0.02 cm/cm (in. /in.)) should be in excess of 1000 cycles. Also, fatigue
life was not very temperature dependent over the temperature range used in these tests.
FABRICATION AND EVALUATION OF BIMETALLIC TUBING AND JOINTS
As mentioned previously, there was two SNAP -8 boiler designs. The bimetal tube
boiler was a counterflow design (fig. 13) using seven Ta/316 SS tubes from one boiler
header to the other. The double containment boiler design (fig. 14) consisted of seven
unalloyed Ta boiler tubes, each within a flattened-oval 321 SS tube, which was required
to accommodate the difference in thermal expansion between the Ta and the 321 SS. The
annulus between each Ta and 321 SS tube was filled with nonflowing NaK. The seven
tubular assemblies were further contained in a 316 SS boiler outer shell within which the
primary loop NaK flowed in a counterflow direction. The Ta mercury containment tubes
were interconnected by Ta header-manifolds at the boiler inlet and outlet. To avoid con-
tinuation of Ta outside the 316 SS boiler outer shell, the Ta header-manifolds were
joined within the shell to 316 SS at the boiler inlet and outlet by means of Ta/316 SS
17
Flow
Flowing NaK
Figure 13. - SNAP-8 bimetallic tubing boiler configuration.
transition joints. Zirconium foil was wrapped around the joints in order to protect them
from embrittlement by any carbon and/or nitrogen that might be in the static NaK.
Both the bimetallic tubing and bimetallic joints required considerable fabrication
process development and evaluation. This was performed under several NASA contracts
and the results are summarized in the following sections.
Initial Investigation
Early in the SNAP -8 program, consideration had been given to the possibility of
using refractory metals such as Nb or Ta to contain the Hg working fluid in the boiler .
It was believed at that time that steel-clad bimetallic tubing would be required to pre-
vent the refractory metal from gettering embrittling elements, such as carbon and ni-
trogen, from the flowing NaK stream. Experimental quantities of Nb or Ta/SS tubing
were produced by several vendors using the following fabrication processes: hot co-
extrusion and drawing, explosion welding and drawing, and explosion welding to size.
A program to evaluate the various typesof bimetallic tubing and to develop welded joints
using this tubing was conducted at the Wes ting house Astronuclear Laboratory (refs. 21
and 22). The various types of tubing were compared on the basis of bond integrity, di-
mensional control, and surface condition. On this basis, the Nb/316 SS tubing produced
by hot coextrusion and drawing was considered to be the best. The weld-joint investiga-
18
(b) Boiler end.
Figure 14. - SNAP-8 double containment boiler configuration.
19
tion was concerned with three basic configurations: a butt joint, a tee joint, and a tube-
to-header joint. Coextruded and drawn Nb/316 SS tubing was used. After considerable
experimentation, a successful design for each of the three basic configurations was de-
veloped (figs. 15 to 17). Several of each type joint were produced and tested primarily
to determine their ability to withstand thermal cycling between 320° and 730° C (600°
and 1350° F) and an internal pressure of 3.9 meganewtons per square meter (565 psia)
at 730° C (1350° F) . The test results indicated that the joints would withstand the
SNAP-8 system operating conditions (370° C (1350° F) and 1.9 MN/m 2 (275 psia)). The
details of the tubing evaluation and joint development efforts are described in references
21 and 22.
Concurrent with these programs, an investigation was conducted to determine the
optimum refractory metal/austenitic alloy combination for use in the SNAP-8 system.
This program was also conducted at the Westinghouse Astronuclear Laboratory (ref. 23).
Sixteen bimetallic combinations, produced by explosion welding, were evaluated by in-
terdiffusion experiments at 760°, 820°, and 870° C (1400°, 1500°, and 1600° F), room
temperature tensile tests, creep-rupture tests at 730° C (1350° F), and thermal cyclic
testing between 320° and 730° C (600° and 1350° F). The bimetals consisted of Nb, Ta,
Nb-lZr, FS-85, or T-222 in combination with 321 and 347 SS or the nickel -base alloys
Inconel 600 and Hastelloy N. A major finding of the program was that the bimetallic
-4
combinations having an interdiffusion zone thickness of less than 12.7x10 centimeter
(5xl0“^ in.) would withstand a minimum of 20 thermal cycles between 320° and 730° C
(600° and 1350° F) without degradation of the interface weld. It was concluded that the
optimum refractory metal/austenitic alloy combination was tantalum/300 series stain-
less steel, primarily on the basis of minimum interdiffusion. The details of the evalua-
tion can be found in reference 23.
Tantalum/316 Stainless Steel Tubing Development
General . - As mentioned previously, evaluation of a small quantity of several simi-
lar types of experimental bimetallic tubing had indicated that the piece of tubing pro-
duced by hot coextrusion followed by cold drawing was the best. Other sections of tubing
from this same lot, however, were found in other investigations to have a significant
amount of nonwelded areas. It was therefore considered necessary to further improve
the fabrication procedures for bimetallic tubing before tantalum/300 series stainless
steel could be seriously considered for SNAP-8 use. Since cold drawing was apparently
responsible for the unwelded areas observed, it was decided to investigate processes
that excluded this step. The processes investigated were hot coextrusion to final size
and explosion welding to final size. The stainless steel selected was type 316. Hot co-
extrusion to size was attempted by two suppliers under NASA contract: the Nuclear
20
21
la) Assembly view.
Figure 16. - Niobium/316 stainless steel bimetalic tubing tee joint (ref. 21).
L Stain less steel internal re-
inforcement bottom section
'-Stainless steel
bottom section
(b) Sequence of welding.
Figure 16. - Concluded.
Metals Division of the Whittaker Corporation and the Metaionics Division of the Kawecki
Chemical Corporation. Nuclear Metals extruded over a tool steel mandrel; Metaionics
used a filled-billet technique. Explosion welding to size was attempted by Aerojet-
Downey, under subcontract to Aerojet-General. The tubing, in each case, was to have
a 1. 65-centimeter (0. 652-in. ) inside diameter and be longer than 4. 6 meters (15 ft).
The Ta wall thickness in all cases was to be 0. 051 centimeter (0. 020 in. ). The 316 SS
wall thickness was to be 0. 152 centimeter (0.060 in. ) for the coextruded tubing and
0.203 centimeter (0.080 in. ) for the explosion welded tubing (readily available stock).
Hot coextruded tubing - mandrel . - The billet configuration for mandrel extrusion
(fig. 18) consisted of four concentric cylinders: carbon steel on the inside, then Ta,
stainless steel, and carbon steel on the outside. Carbon steel front and rear plates and
Low carbon steel-
y-Copper plating
Evacuation tube-'
Low carbon steel
^Copper plating
^ Low carbon steel
— Low carbon steel
(normally tack
welded to main
billet)
l Low carbon steel
l 316 Stainless steel
Figure 18. - Extrusion billet design for producing tantalum/316 stainless steel tubing using a mandrel extrusion method.
a carbon steel nosepiece were also used. The assembly was welded as shown, then
outgassed and sealed at an elevated temperature . The original overall billet size was
calculated to produce a tube about 6. 1 meters (~20 ft) long at a reduction ratio of 32:1.
The final extrusion conditions selected were 1000° C (1830° F) and a reduction ratio of
25:1. The particulars of the billet assembly process and the extensive development
program required to produce sound tubing, heretofore unpublished, are presented in the
appendix. Thirteen tubes, approximately 4.6 meters (15 ft) long, were successfully
produced and prepared for subsequent evaluation (to be discussed in a later section) .
Hot coextruded tubing - filled billet. - The filled-billet configuration (fig. 19) con-
sisted of a plain carbon steel core over which was tightly fit, in turn, a Ta cylinder and
a stainless-steel cylinder . A nose piece and tail piece, also of carbon steel, were
24
Figure 19. - Extrusion billet design for producing tantalum/ 316 stainless steel tubing using a filled billet method.
welded to the 316 SS outer cy liner . The assembly was then outgassed and sealed at an
elevated temperature. The overall billet size was calculated to produce a tube about
6.1 meters (20 ft) long at a reduction ratio of 20:1. The nominal extrusion temperature
was about 1030° C (1880° F). The particulars of the billet assembly process and the ex-
trusion process, heretofore unpublished, are presented in the appendix. Twenty tubes,
greater than 4.6 meters (15 ft) long, were successfully produced and prepared for sub-
sequent evaluation (to be discussed in a later section) .
Explosion welded tubing . - The billet configuration used for explosion welding is
shown in figure 20. The Ta tubing was dimpled by means of a special hydraulic tool.
After degreasing, the inside of each tube was filled with Cerrobend-A using a vertical
casting technique. Each tube was inspected to insure a void-free Cerrobend core. The
bonding was performed with the Ta and 316 SS tubing fixed vertically in a hole in the
ground. The explosive used was nitroguanadine powder packed around the outside diam-
eter of the 316 SS tube and held in place by a 8. 9 -centimeter (3. 5-in.) diameter sur-
rounding cardboard tube. The explosion welding technique was successfully developed
to the point that 4.6-meter (15 -ft) long tubes could be produced. The process, however,
had an inherent, undesirable characteristic, which was the lack of welding at the dimple
locations. A more detailed description of the explosion welding procedure can be found
in reference 24. Thermal cyclic testing of both the explosion welded tubing and the
coextruded tubing is described in reference 1.
Evaluation of tantalum/316 stainless steel bimetallic tubing . - Specimens of all
three types of coextruded tubing were thoroughly evaluated at Westinghouse Astronuclear
Laboratory under NASA contract (ref. 24). All of the tubing was carefully dimensionally
inspected: outside diameter, inside diameter, length, and straightness. In addition,
each tube was inspected by dye penetrant and ultrasonic techniques.
25
Typical dimple,
0.318 cm (0.125 in. )
diam by 0. 05 cm
(0.020 in.) high —
Area of contact
is not bonded
in final bimetal
Tube filled with
cerrobend alloy
Standoff
-Tantalum tube
-316 Stainless
steel tube
-Cavity filled with
expolosive powder -
detonation cap at
one end
Cardboard
outer
container
Figure 20. - Configuration and setup for explosion welding of tantalum/316 stainless
steel bimetallic tubing.
All three types of tubing had good dimensional control (figs. 21 to 23), although the
inside diameter of the coextruded filled -billet tubing was extremely rough (fig. 21).
The dye penetrant inspection revealed numerous inside and outside diameter surface
defects on much of the coextruded filled-billet tubing . But the mandrel extruded and
explosion welded tubing showed no surface defects. Ultrasonic inspection showed the
filled-billet tubing to have a few nonwelded spots. And, as expected, the explosion
welded tubing was nonwelded at the dimple locations, again, as determined by ultrasonic
inspection. The mandrel -extruded tubing had no nonwelded defects. Metallography re-
vealed that both of the coextruded types of tubing had continuous welds and no significant
intermetallic layers were observed at the bimetal interfaces. Conversely, the explo-
sion welded tubing showed a nearly continuous hard intermetallic layer at the bimetal
interface resulting from the incipient melting that occurs during the explosion welding
of the two surfaces .
The test program consisted of thermal cycling between 320° and 730° C (600° and
1350° F) for 100 cycles to determine the effect of thermal stress on the bimetal weld
interfaces. Biaxial creep burst tests at 730° C (1350° F) were also conducted using in-
ternal gas pressurization. It was observed that the nonwelded spots on the explosion
welded tubing propagated along the weld interface during thermal cycling. The coex-
26
Figure 21. - Transverse section of Tantalum/316 stainless
steel tubing produced by a filled billet method (ref. 24).
Un etched.
Figure 22. - Transverse section of Tantalum/316 stainless steel tubing pro-
duced by a Mandrel extrusion method (ref. 24). Unetched.
Figure 23. - Transverse section of Tan.talum/316 stainless
steel tubing produced by explosion welding (ref. 24).
Unetched.
truded types showed no noticeable change. After considering all test results for the
three types of tubing, it was concluded that the tubing coextruded over a mandrel was the
best primarily because of the combination of good surface finish, dimensional control,
and general integrity during testing. The details of the evaluation program can be found
in reference 24.
Tantalum/316 Stainless Steel Transition Joint Development
General . - As stated previously, the double containment boiler design required
Ta/316 SS inlet and outlet transition joints. To satisfy initial requirements, trial joints
of this type were procured from two suppliers: the General Electric Company and the
Nuclear Metals Division of the Whittaker Corporation. The General Electric joints were
produced using a brazed tongue- in-groove design. The braze alloy used was J-8400
(Co-21Cr-21Ni-8Si-2. 5W-0. 8B-0. 4C). The Nuclear Metals joints were hot coextruded
over a mandrel and had a tandem, tapered interface design. Several joints of each type
were produced and evaluated. Both types of joints were operated in three full-scale
SNAP-8 boilers that were built at Lewis. Also, testing of several joints of each type
was performed by Aerojet General Corporation. Results of these tests indicated that
both joint types required additional fabrication optimization before being considered
adequate for 100 startup cycles and 5 years of service at 730° C (1350° F). During the
Aerojet testing, some of the brazed joints showed a serious lack of braze filling. Some
of the coextruded tandem joints showed flaws at the interface between the Ta and 316 SS
on the outer circumference, which propagated during test. Asa result of these deter-
minations, programs were initiated to optimize the fabrication procedure for each joint.
28
A third joint design, best described as a sleeve joint, was added to this development
effort. Basically, the new joint was a large diameter, heavy -wall bimetallic tube ma-
chined to a joint configuration, and thus it was basically an extension of the technology
acquired during the development and testing of the coextruded bimetallic tubing. It of-
fered several potential advantages over the other types of joint. For example, the sleeve
joint, as a bimetallic tube, could be extended toward the SNAP-8 turbine or Hg pump if
corrosion effects in the Hg vapor line to the turbine, or liquid Hg line from the pump,
became significant. This extension was, of course, not possible with the other two
joints. Another advantage would be the long interface length. If interfacial separation
occurred, it would have to propagate a much longer distance with the sleeve joint than
with the tandem (3. 81 cm (1,50 in.) long interface) or brazed joint (1. 90 cm (0. 75 in.) long
interface) . Also, a sleeve joint would be much easier to inspect by ultrasonic techniques
than either of the other two. Another important advantage would be that a piece of the
actual joint could be destructively , as well as nondestructively, examined in the as-
fabricated condition by merely removing a ring from either end. It was obviously not
possible to destructively examine either the tandem or brazed joint in the as-fabricated
condition.
The brazed joint fabrication optimization program was conducted by the General
Electric Company. The fabrication optimization program for the two coextruded joints
was conducted by Nuclear Metals. The three types of joint and their dimensions are
shown in figure 24.
Brazed joints . - A major change to the brazed joint configuration was made for this
program. In the early brazed joint design, the tongue of the tongue -in -groove design was
stainless steel and the groove was Ta.. For the joints produced under this program, this
was reversed: the Ta became the tongue and the stainless steel the groove. This ar-
rangement was considered more satisfactory from several standpoints, which included
lower joint stress, better braze filling, and improved cleaning efficiency before brazing.
The brazing process consisted of several steps. The components were assembled in a
vertical position, and the braze alloy (J-8400) was applied as a slurry. The assembly
was then positioned in a vacuum furnace, and the furnace was evacuated to a pressure of
less than 6.7x10"° newton per square meter (0. 5x10 ° torr). The assembly was heated
to the brazing temperature, held at temperature for a brief time, and then cooled slowly.
Extensive experimentation revealed that minimum braze microshrinkage could be
achieved by brazing either at 1180° C (2160° F) for 5 minutes or at 1230° C (2250° F)
for 1 minute and cooling at a rate of 15° C (25° F) per minute during braze solidification.
Another program goal, in addition to determining an optimum fabrication procedure,
was to ascertain a reliable ultrasonic inspection method for determining the quality of
the joints. The capability of ultrasonics to accurately delineate braze integrity was
demonstrated by correlating inspection data with physical microstructures of actual pro-
totype joints.
29
Tantalum
4.47
U. 760)
i.d.
Brazed Joint
y— 316 Stainless steel
Tandem tapered joint
Electron-
beam weld-\ ,f - 316 Stainless steel
Figure 24. - Tantalum/316 stainless steel transition joint configuration.
(All dimensions are in cm (in. ). )
Figure 25. - Longitudinal section of brazed joint in the as-brazed condition (ref. 27).
Twelve 5. 1 -centimeter (2 -in.) outside diameter tubular joints were successfully
brazed, and their quality verified by ultrasonic inspection. A typical brazed joint longi-
tudinal cross section is shown in figure 25. The details of this fabrication optimization
program can be found in reference 25.
Hot coextruded joints . - The extrusion optimization program for the coextruded tan-
dem and sleeve joints included scrupulous control of prebillet assembly cleaning proce-
dures and careful outgassing and sealing procedures. Also, Ta foil was used in the billet
assembly to getter any entrapped air. Several trial billets of each type of joint were ex-
truded at different temperatures and extrusion ratios. Based on the examination of the
trial extrusions, the extrusion conditions for the final joints were decided upon. The
tandem joints were extruded at 995° C (1825° F) at a 5:1 extrusion ratio; the sleeve
joints were extruded at 1070° C (1950° F) at an 8:1 extrusion ratio. Twelve tandem
joints were produced. A longitudinal cross section of a typical joint is shown in fig-
ure 26. Eight sleeve extrusions were made from which 16 sleeve joints were obtained,
since each extrusion was over 61.0 centimeters (24 in.) long. The sleeve cross section
was similar to that for coextruded tubing (fig . 22) except that the total wail thickness was
0.89 centimeter (0.35 in.).
Figure 26. - Longitudinal section of coextruded joint in the as-extruded condition
(ref. 27). Unetched.
31
In addition, two smaller sleeve extrusions were produced with 1.91 -centime ter
(0.75-in.) inside diameters and the same Ta and 316 SS thicknesses as the large sleeve
extrusions. Because these extrusions were over 1.0 meter (40 in.) long, three sleeve
joints were obtained from each. These joints were designed to be used at the Hg inlet of
the double containment boiler. The details of this fabrication optimization program can
be found in reference 26 . *
Evaluation of joints . - Several joints of each of the three joint types were tested and
thoroughly evaluated by Westinghouse Astronuclear Laboratory under Lewis contract
(ref. 27). In order to simulate the stress imposed by the double -containment boiler end
flange, a stainless-steel collar was attached to the sleeve joints by means of electron
beam welding. Each joint was carefully dimensionally inspected (outside and inside di-
ameters, length, and straightness). In addition, each joint was inspected by helium leak,
dye penetrant, and ultrasonic techniques. The wall thickness of the sleeve joints varied
considerably because the as -extruded sleeves had been insufficiently straightened before
machining. The taper lengths of the tandem joints varied from 2. 56 to 4.24 centimeters
(1.01 to 1.67 in.). The dimensions of the brazed joints showed no significant variation.
A few minor surface defects were discovered in each group of joints by the dye penetrant
technique. But helium leak checking and the ultrasonic inspection showed no significant
potential problem areas .
The joint testing program consisted of thermal cycling four of each type of joint be-
tween 120° and 730° C (250° and 1350° F); two of each type of joint were unpressurized,
and two were pressurized to 1.83 meganewtons per square meter (265 psia) with argon
to simulate boiler operation. Each joint was cycled 100 times; the holding times at
730° C (1350° F) varied from 2 to 10 hours, but two 100-hour soaks were included. In
addition, several of the joints were subjected to a 1000-hour soak at 730° C (1350° F)
prior to their final 10 cycles. The joints were thoroughly nondestructively inspected
periodically by the aforementioned techniques as the thermal cycling progressed.
None of the joints tested showed leaks as determined by the helium leak check. But
the dye penetrant and ultrasonic inspection revealed that the condition of all joints de-
teriorated somewhat during testing. The coextruded sleeve and tandem joints were con-
siderably worse than the brazed joints in this regard. The sleeve and tandem joint de-
terioration consisted primarily of fissures at or near the Ta/316 SS interface. The
brazed joints displayed a small amount of microcracking in the braze. An illustration
of the fissuring is shown in figure 27. In addition, the sleeve joints bowed somewhat
during test, and both the sleeve and tandem joints displayed diametral contraction in the
Also in this program, an attempt was made to produce 1.65-cm (0.652-in.) i.d.
bimetallic tubing using thin, rolled -up fine-grained Ta sheet in place of seamless Ta
tubing and a filled -billet technique. This effort was not successful.
32
<a> inside-diameter fissure in stainless steel area.
(b) Outside-diameter fissure in Tantalum area.
Figure 27. - Longitudinal sections of coextruded joint after thermal cycling, showing bond line fissures.
bimetal transition area. No significant interdiffusion between the Ta and 316 SS or Ta,
braze, and 316 SS was observed as a result of temperature exposure, and none of any
significance would be expected over a 5-year period, based on calculated diffusion rates.
It was concluded that, although none of the joints actually failed in test, the sleeve and
tandem joints indicated a need for future development before they could be considered
acceptable for SNAP-8 service. The brazed joint, having demonstrated good dimen-
sional stability and joint durability was concluded to be the best candidate for SNAP -8
use (at the current state of development) .
This conclusion was reinforced by a review of some earlier testing of similar
Ta/316 SS brazed joints. A 10-centimeter (2.5-in.) outside diameter brazed joint had
33
survived 150 severe thermal shocks in a high flow velocity Hg test system without ap-
parent damage (ref. 28). Also, failures during 730° C (1350° F) tensile tests of brazed
joint configurations had never occurred in the braze itself (ref. 17). Finally, several
brazed joints had been used in Hg boiler tests with good results from both a corrosion
and structural standpoint.
APPLICABILITY OF RESULTS
Although work on the SNAP-8 pregram has been terminated, it should be recognized
that much of the materials technology that was developed could well be applicable to
other systems. For example, a mercury -rankine system is being developed to power an
artificial heart. Investigators involved in this development are using corrosion informa-
tion generated under the SNAP-8 program. Also, it is likely that some of the NaK cor-
rosion information could be used in nuclear reactor systems for land-based power (e.g. ,
breeder reactors).
The Ta tensile, creep, and low-cycle fatigue testing has thoroughly described this
material’s properties in a temperature range where previously little data were avail-
able. Also, basic designs for refractory metal /stainless steel bimetallic welds were
developed; these could easily be applied to other dissimilar bimetallic systems. Var-
ious types of refractory metal/stainless steel bimetallic tubing produced by techniques
developed under the SNAP-8 program could be used in heat-pipe applications. Or, with
a reduced amount of refractory metal, it could be used in commercial chemical systems
where the corrosion resistance of refractory metals would extend tubing life greatly.
Refractory metal /stainless steel bimetallic transition joints produced by techniques de-
veloped under the SNAP-8 program are being utilized in thermoelectric modules for
space power systems and could be used in other advanced power systems. Many other
applications of the technology developed are, of course, possible.
SUMMARY OF RESULTS
Several major conclusions can be drawn based on the materials technology efforts
summarized in this report:
1. Tantalum, niobium - 1-percent zirconium, and alloy T- 111 (Ta-8W-2Hf), unlike
more conventional cobalt and iron-base alloy containment materials such as L-605 and
9 chromium-1 molybdenum steel, should not be affected by liquid Hg exposure at tem-
peratures at least up to 650° C (1200° F).
2. The sodium -potassium eutectic alloy (NaK) corrosion rates of the major SNAP -8
primary reactor loop materials Hastelloy N and 316 SS appear to be acceptably low (less
34
than 0.004 cm/10^ hr or 0.0015 in. /10^ hr) at temperatures up to 700° C (1300° F) pro-
viding the oxygen level in the NaK is maintained at less than 30 ppm.
3. Tantalum having a uniformly distributed oxygen concentration of 115 ppm or less
will not be attacked by NaK at temperatures up to 730° C (1350° F) .
4. Uniaxial creep testing of Ta tubing at 730° C (1350° F) revealed a strong depen-
dence of creep strength on grain size, the fine-grained tantalum being considerably more
creep resistant.
5. Low-cycle fatigue testing of unalloyed Ta bar at temperatures up to 730° C
(1350° F) revealed that the Ta low-cycle fatigue life at the maximum plastic strain range
(0.02 cm/cm (in. /in.)) to which it would be exposed in the SNAP-8 system should be in
excess of 1000 cycles at temperatures up to 730° C (1350° F) .
6. Several different welded joint designs using refractory metal /austenitic alloy bi-
metallic tubing can be successfully produced. The three basic configurations produced
in this program were a straight butt joint, a tee joint, and a tube-to-header joint.
7. The most suitable refractory metal/austenitic alloy bimetallic couple for fabri-
cation into tubing for mercury containment service at temperatures up to 730° C
(1350° F) was determined to be tantalum/type 316 stainless steel (Ta/316 SS).
8. The preferred fabrication method for producing Ta/316 SS bimetallic tubing was
determined to be hot coextrusion over a mandrel. The two other techniques attempted,
hot filled -billet coextrusion and explosion welding, had significant deficiencies.
9. A brazed Ta/316 SS tubular bimetallic transition joint is considered to be the best
candidate for SNAP-8 use at the current state of development. The other two types of
joint, the hot -coextruded sleeve and tandem, require further development before they
could be considered acceptable for SNAP -8 service.
Lewis Research Center,
National Aeronautics and Space Administration,
Cleveland, Ohio, May 7, 1973,
501-21.
35
APPENDIX - PROCEDURES USED TO COEXTRUDE TANTALUM/316
STAINLESS STEEL B I METALL I C TUB I NG
In order to provide the SNAP- 8 system with reliable Ta/316 SS bimetallic tubing for
use in the mercury boiler, it was necessary to optimize tubing fabrication techniques.
A program was conducted by the Nuclear Metals Division of the Whittaker Corporation
that utilized hot extrusion over a mandrel to produce the tubing, and it involved a fairly
extensive study of the extrusion variables associated with this process. The other pro-
gram, conducted by the Metaionics Division of the Kawecki Chemical Corporation, in-
volved hot extrusion of a filled billet. This study was much less extensive and was
basically limited to applying extrusion conditions previously developed for other applica-
tions. The purpose of this appendix is to describe these two extrusion programs since
neither of them have been reported previously.
MANDREL EXTRUS ION TECHNIQUE
Billet Component Preparation
After machining to the sizes indicated in figure 28, the seamless mild -steel compo-
nents were outgassed at about 1010° C (1850° F) for 24 hours in a heated vacuum retort.
Following cooling and removal from the retort, they were stored in plastic bags with a
dehumidifying agent until required for assembly.
Similarly, the machined 316 SS sleeves were scrubbed with a detergent solution,
and rinsed with tap water, distilled water, acetone, and ethanol.
The Ta sleeve was machined to the size needed for the first extrusions. It was de-
greased with trichloroethylene and then acetone. Then it was chemically etched in a
solution composed of one volume hydrofluoric acid (49 percent assay), two volumes
sulfuric acid (96 percent assay), and two volumes nitric acid (70 percent assay). This
solution was used to remove a 0.0025- to 0. 0051-centimeter (0.001- to 0.002-in.) thick
surface layer from the Ta. The chemical etch was followed by a thorough tap water
rinse, a distilled water rinse, an acetone rinse, and an alcohol rinse.
Immediately after the Ta was cleaned, the billets were assembled into the configura-
tion shown in figure 28. First the steel end closures (one with an evacuation tube) were
welded, and then the shaped nose piece was attached. After helium leak checking, the
billet was ready for copper plating, outgassing, and sealoff.
The assembled billets were electroplated with copper by a standard procedure that
began with a thorough degreasing of the exterior in a trichloroethylene bath. The clean
36
,635 (0.25)
/-Low carbon steel
.✓'-Tantalum
✓—Low carbon steel
/- Copper, 0. 013 (0. 005) thick
—Low carbon steel
— Low carbon steel
(normally tack
welded to mai n
billet)
\ u 316 Stainless steel
'-Evacuation tube
figure 28. - Extrusion billet design for producing tantalum/316 stainless steel tubing using a mandrel extrusion method. Extrusion ratio, 25:1.
(All dimensions are in cm (in. ). )
billet was then dipped in a hydrochloric acid solution, rinsed in tap water, and trans-
ferred to a copper cyanide plating bath where it received a flash coating of copper. The
final coating of copper, 0.038 centimeter (0.015 in.) thick on the outer surface of the
tubular billet and 0.0076 to 0.0127 centimeter (0.003 to 0.005 in.) thick on the interior
of the billet, was applied in an acidified copper sulfate bath. The evacuation tube was
protected during plating by a wrapping of electrical tape.
-2
The billets were attached to a vacuum system and evacuated to 1.33x10 newton
per square meter (10"^ torr) and then slowly heated to 430° C (800° F). After at least
4 hours at 430° C (800° F) or until outgassing was completed, the billets were slowly
cooled to room temperature . Sealoff was accomplished by torch heating the evacuation
tube to bright red heat, then hammering it flat close to the end plate, and melting off
the excess. The copper on the sealed -off billets was abraded to remove oxide formed
during outgassing, washed with trichloroethylene , then spray -coated with a dry graphite
film lubricant, after which the billets were ready for loading into the furnace.
Extrusion Procedures
The tools, 18-4-1 steel (hardened to Rockwell C 58 to 60 for the backers and man-
drels) and H-21 or M-2 steel for the dies, were degreased with trichloroethylene, then
heated in 480° C (900° F) furnaces for several hours until coated with an oxide film.
37
They were then cooled and sprayed with a graphite lubricant. In addition, the dies were
coated with Necrolene. The liner was wire brushed and blown free of any debris.
The billets were heated in stainless steel retorts flooded with flowing argon. The
retorts containing the billets were heated for a minimum of 3 ^ hours in a resistance
furnace.
Extrusions were made in a 1.2 5 -meganewton (1400 -ton) hydropress, using a 7.72-
centimeter (3. 050 -in.) inside diameter liner. The maximum force available to the liner
was 6.8 meganewtons (770 tons).
Overheating and softening of mandrels was minimized by attaching them to the stem,
which was attached to the main ram of the press. A roll pin was generally adequate to
maintain proper alinement of the mandrel. Further assurance of proper alinement was
provided by a graphite sleeve, which tightly fit both the stem and the mandrel backer.
The sequence of operations was as follows:
(1) Attach mandrel to stem.
(2) Insert die in liner 7 to 10 minutes before extruding.
(3) Lubricate mandrel and liner.
(4) Remove billet from furnace and hand load until nose enters liner.
(5) Extrude at a reduction ratio of 32:1.
Several tubes were extruded over the temperature range 940° to 1090° C (1725° to
2000° F) . Extrusion speeds varied over the range 2. 5 to 12.7 meters per minute (100 to
500 in. /min) . A basic process was demonstrated in that tubing was successfully ex-
truded nearly to the specified dimensions. Based on the tube surface condition, the op-
timum extrusion temperature was determined to be 995° C (1825° F). The extrusion
speed selected was 7.6 to 12.7 meters per minute (300 to 500 in. /min).
initial Results
Samples of the tubing subjected to microscopic study generally revealed a bondline
layer less than 5xl0~ 5 centimeter (2xl0~ 5 in.) thick, but with protuberances as thick as
1x10"^ centimeter (4x10 in.).
When split lenghtwise, 5.1 -centimeter (2-in.) long samples removed from the mid-
dles and the ends of several tubes revealed that the Ta surface was smooth and appar-
ently defect free. A more complete examination of these and later tubes with a bore-
scope revealed surface defects that appeared to be incipient tears. These appeared to
be random in location in the tube, but they were alined longitudinally quite frequently.
Although these tears appeared to be only about 0.005 centimeter (0.002 in.) deep, they
were considered to be symptomatic of a potentially serious deficiency in the basic pro-
cess. A detailed investigation into the causes of these tears was therefore conducted.
38
Process Improvement
The random occurrence of relatively localized defects suggested some defect in the
materials composing the billet rather than in the extrusion technique . The Ta, which
was rather large grained , was especially suspect. Unfortunately , all of the Ta stock
was procured at the start of the program, hence, that parameter could not be varied.
The tears could, however, also have been attributed to many other process variables
and these were investigated.
Honing of the Ta and heavy etching was first tried, but this produced no improve-
ment. Then, the basic mismatch in stiffness of the components was checked by extrud-
ing sheathed rods of Ta in tandem with stainless steel and comparing the extrusion con-
stants. No serious mismatch problems were apparent between the materials used in the
first billets; moreover, it did not appear possible to obtain a better match by adjusting
the temperature within the range of extrusion temperatures available.
Stiffness of the inner steel extrusion sheath was investigated by varying three pa-
rameters. The first change from the basic process was to increase the thickness of the
sheath to offset chilling from contact with the cold mandrel and to thereby provide less
of a stiffness gradient through the sheath to the tantalum. Thick sheathing also provided
assurance that the sheath was not tearing because of being extruded too thin. Secondly,
a warm mandrel was used to further counteract chilling of the sheath. A third variation
involved changing to a sheath material with a different stiffness than the low carbon steel.
The first trial extrusion showed that roughly double the inner sheath thickness was
not beneficial. The next extrusion indicated that a double sheath thickness with a warm
mandrel might be a step in the right direction. Continued improvement of the inside -
diameter surface of succeeding extrusions occurred as the inner sheath was increased in
thickness. Other attempts to adjust the stiffness of the inner sheath by using materials
other than low carbon steel or by increasing the thickness of the copper outside of the
steel resulted in no marked improvement. Unsuccessfully used as inner sheath materials
were copper, copper - 30-percent nickel, and Monel .
Lowering the extrusion reduction, which frequently alleviates tearing problems, was
explored as another alternative. This appeared to result in a better Ta surface.
Other changes were also incorporated into the process. The first billets were
heated on their sides in argon -flooded stainless steel retorts. In this position, forced
contact of billet components might permit localized interaction of the components, and
this could conceivably produce embrittlement of the Ta surface , which would then tear
while being extruded. Opportunity for such interaction was minimized by heating sub-
sequent billets vertically (standing on their noses) either inside graphite cans or directly
exposed to the nitrogen furnace atmosphere.
All but one vertically heated tube appeared to be tear-free. Therefore, although
the conclusion that vertical heating prevented tears could not be drawn, it was decided
39
to include it in the process. Heating inside graphite cans or directly in nitrogen did not
appear to make a difference in the quality of the extrusions.
The results of this study led to the following conclusions:
(1) Changing temperatures in the range 940° to 1090° C (1725° to 2000° F) with rela-
tively thin inner steel extrusion sheaths did not appear to prevent tears in the Ta.
(2) Softer (all copper) or staffer (Monel) inner extrusion sheaths did not produce a
good Ta surface.
(3) The effect of various inner sheathing materials merits further investigation, al-
though both copper (a soft material) and Monel (a stiff material) permitted surface tears.
One possibility is the use of 316 SS for the inner sheath. This could be readily dissolved
by aqua regia without affecting the external 316 SS.
(4) The use of a warm mandrel in combination with a low reduction ratio and a
thick steel inner sleeve (0.51 cm (0.2 in.) wall) minimized or eliminated tearing.
Postextrusion Processing
Some of the early tubes were stretch -straightened with less than 1.5 percent per-
manent strain. The straightening was done on a hydraulic draw bench equipped with two
sets of grips, one on the ram and one attached to the rear of the bench. The ends of the
tube were fitted with steel plugs to prevent the tube from collapsing. It was found that a
1.2- to 1. 5-percent permanent strain was adequate to straighten the tubes.
Removal of the bulk of the extrusion jacket was accomplished by use of a nitric acid
solution. Concentrations of 30 to 50 volume percent of nitric acid (70 percent assay)
with water were suitable. Frequently, pinhead-size pieces of steel became passivated.
These were attacked by numerous methods such as heating the nitric acid solution to
boiling. Hydrochloric acid appeared to be capable of removing the traces of steel if
given enough time. Aqua regia seemed to perform most rapidly; however, there was
some indication that the steel could also become passivated to this. All of the methods
for removing the steel were intermingled for most of the tubing made in this pregram.
The tubes were subjected to successive treatments until nothing that could be suspected
of being steel could be seen on the Ta. Further study of techniques for the removal of
passivated steel in such tubes is desirable.
A final chemical polish of the Ta was obtained using a solution of the same composi-
tion as that used to prepare the Ta billet components for extrusion. About 10 minutes in
this solution at 40° C (110° F) cleaned the tubes and appeared to remove a layer of Ta
about 0.0025 centimeter (0.001 in.) thick. This procedure also needs improvement.
More uniform attack would probably occur if the solution were slowly circulated through
the tube by an air lift or by an acid pump. During the program the tubes were rolled
and occasionally partially drained and then refilled by raising and lowering their ends.
40
The chemical polish was completed by transferring the tubes to a cold water rinse,
followed by a hot water rinse . The tubes were then raised to a vertical position to drain .
Acetone poured through the tube hastened the drying.
The straightened tubes were centerless polished on an abrasive -belt, centerless
polishing machine. A 10. 2 -centime ter (4-in.) wide, 320-grit silicon carbide belt was
adequate to remove about a 0.0025- to 0.0076 -centimeter (0.001- to 0.003-in.) thick
surface layer in one pass and leave a 16 rms surface finish. A water based coolant was
used. A 4. 6 -meter (15-ft) tube required about 10 minutes to pass through the machine.
Final Processing Conditions
The development program resulted in procedures that could produce straight tubing
with the tantalum - stainless steel well welded and with polished stainless steel sur-
faces. The surface of the Ta was not as smooth as desired, but no tears were ob-
servable. Since both a heavy inner steel extrusion sheath and a lower reduction pro-
duced tear-free Ta, these two techniques were combined to form the basis of the subse-
quent processing.
In order to verify the final process, a trial batch of three tubes was made and the
interiors were carefully examined for tears. About 3.7 meters (12 ft) of one tube was
split lengthwise to preclude any confusion due to bore sc ope inspection. All tubes were
tear -free; some heavy ripples were apparent near the ends of two of the tubes and near
the middle of the third tube .
The quality of the three tubes was considered an adequate verification of the extru-
sion process so that the remaining stock could be extruded. The ripples in the Ta sur-
face indicated that the process was not, however, fully optimized.
The billets for the final tubing utilized a 0. 51-centimeter (0. 2-in. ) thick inner steel
extrusion sheath. The billets were extruded at 995° C (1825° F) at a reduction ratio of
25:1 and an extrusion speed of 12. 7 meters per minute (500 in. /min). The tool steel
mandrels were heated to 480° C (900° F).
After straightening, removal of the extrusion jacket by pickling, and chemical and
mechanical polishing, the tubes were given a final inspection and found to be acceptable.
FILLED-B ILLET EXTRUSION TECHNIQUE
All components were machined to the billet design shown in figure 29. No ma-
chining of the outside diameter of the stainless was necessary. Following all the ma-
chining operations, each steel component was vapor degreased with trichloroethylene
and then degassed at 1065° C (1950° F) in a vacuum of 1.33x10“^ newton per square
41
Figure 29. - Extrusion billet design for producing tantalum/316 stainless steel tubing using a filled billet method.
(All dimensions are in cm (in. ). )
meter (10”^ torr). After the vacuum degassing operation, each part was cleaned in ace-
tone and assembled. Careful attention was paid not to get contaminants on any of the
pieces before assembly. The unit was then welded together. Sealing was performed at
540° C (1000° F) in a vacuum of 1. 33x10"^ newton per square meter (10“® torr).
After evacuation and sealing, the billets were inserted into a furnace which was at
the prescribed extrusion temperature. The furnace door was sealed, and argon was
pumped into the furnace. The heating time was approximately hours. Graphite cut-
offs were inserted into the furnace with the billets. The billets were then extruded at
1.0 meter per minute (40 in. /min) through zirconium oxide coated die (2.1 cm
(0.0830 in.) diam) at a 45° approach angle. The force required to start the push was
5. 9 meganewtons (660 tons) rising to a peak of 6. 8 meganewtons (770 tons). A9.1-
centimeter (3. 600-in. ) liner, and a graphite lubricant were used. The billet extrusion
temperatures ranged between 990° and 1060° C (1810° and 1940° F) although most were
extruded at 1030° C (1880° F). Use of the higher extrusion temperature appeared to
result in an overall better surface finish.
After extrusion, the rods were straightened while they were still hot and allowed to
cool in air to room temperature. The ends were trimmed to leave approximately
4.9-meter (16 -ft) lengths, and the cores were removed. Core removal was accom-
plished by nitric acid leaching at approximately 90° C (200° F) . The acid was pumped
into the tubes.
42
The outside diameter was belt ground to a 2.07-centimeter (0. 815-in.) diameter fol-
lowing a cold hand straightening operation. The inside diameter was sized by pulling a
torpedo type of mandrel through the tube. About four to five passes were necessary to
achieve the size required. Only protruding hills of Ta (0.0127 to 0.0203 cm (0.005 to
0.008 in.)) on the inside diameter were displaced. The tubes were thoroughly vapor de-
greased following the sizing operations.
The oxide formed in the core removal operation was removed from the inside sur-
face by a honing operation. Specifically, 80 -mesh aluminum oxide was blown through the
tube using tank nitrogen.
43
REFERENCES
1. Anon.: SNAP-8 Electrical Generating System Development Program. Aerojet-
General Corp. (NASA CR-72860), July 15, 1971.
2. Rosenblum, Louis; Scheuermann, Coulson; Barrett, Charles A. ; and Lowdermilk,
Warren H . : Mechanism and Kinetics of Corrosion of Selected Iron and Cobalt
Alloys in Refluxing Mercury. NASA TN D-4450, 1968.
3. Vary, Alex; Scheuermann, Coulson M. ; Rosenblum, Louis; and Lowdermilk,
Warren H. : Corrosion in a Cobalt Alloy, Two-Phase Mercury Loop. NASA TN
D-5326, 1969.
4. Owens, James J. ; Nejedlik, James F. ; and Vogt, J. William: The SNAP-2 Power
Conversion System Topical Report, No. 7. Rep. ER-4103, TRW, Inc. , Oct. 26,
1960.
5. Nejedlik, James F. : The SNAP-2 Power Conversion System Topical Report No. 14.
Rep. ER-4461, TRW, Inc., 1962.
6. Cooper, D. B. ; and Vargo, E. J. : Operation of a Forced Circulation, Croloy 9M,
Mercury Loop to Study Corrosion Product Separation Techniques. NASA CR-217,
1965.
7. Farwell, B. E.; Yee, D.; and Nakazato, S. : A 9Cr-lMo Steel as a Mercury Con-
tainment Material for the SNAP -8 Boiler, Rep. 3661, Aerojet -General Corp.
(NASA CR-72503) , Jan. 1968.
8. Weeks, J. R. : Liquidus Curves and Corrosion of Fe, Cr, Ni, Co, V, Cb, Ta, Ti,
and Zr in 500° - 750° C Mercury. Corrosion, vol. 23, no. 4, Apr. 1967,
pp. 98-106.
9. Engle, L. B. , Jr,; and Harrison, R. W. : Corrosion Resistance of Tantalum,
T-lll, and Cb-lZr to Mercury at 1200° F. NASA CR-1811, 1971.
10. Derow, H. ; et al. : Evaluation of Tantalum for Mercury Containment in the SNAP -8
Boiler. Rep. 3680, Aerojet-General Corp. (NASA CR-72651), Nov. 1969.
11. Chalpin, E. S. ; and Lombard, G. L. : Operation and Post-Test Inspection of the
SNAP-8 Pre -Prototyped Boiler P/N CF 751840, S/N2. Rep. AGC-TM-4967:
70-616, Aerojet-General Corp. (NASA CR-72960) , Feb. 24, 1970.
12. Hendrixson, W. H. ; and Harrison, R. W, : SNAP- 8 Refractory Boiler Develop-
ment. Topical Report 3: Evaluation of SNAP -8, SN-1 Boiler. Rep. GESP-550,
General Electric Co. (NASA CR-72814), Nov. 3, 1970.
44
13. Young, Stanley G. ; and Johnston, James R. : Accelerated Cavitation Damage of
Steels and Superalloys in Liquid Metals. NASA TN D-3426, 1966.
14. Kemppainen, D. J. ; and Hammitt, F. G. : Cavitation-Erosion Characteristics of
Selected Materials in Mercury at 500° F. Rep. TR-03424-21-T, Univ. Michigan
(NASA CR- 87000) , Apr. 1967.
15. Savage, H. W.;etal. : SNAP -8 Corrosion Program. Rep. ORNL-3898, Oak Ridge
National Lab. (NASA CR-69822), Dec. 1965.
16. Harrison, R. W.: SNAP -8 Refractory Boiler Development: Corrosion of Oxygen
Contaminated Tantalum in NaK . NASA CR- 1.850 , 1971.
17. Spagnuolo, Adolph C.: Evaluation of Tantalum -to-Stainless -Steel Transition Joints.
NASA TMX-1540, 1968.
18. Spagnuolo, Adolph C.; and Stone, Phillip L, : Tensile Properties of Unalloyed Tan-
talum at Temperatures Up to 1350° F (1000° K). NASA TM X-52670, 1969.
19. Sheffler, K. D. : Generation of Long Time Creep Data on Refractory Alloys at Ele-
vated Temperatures. Rep. TRW-ER-7442, TRW, Inc., 1970.
20. LaForce, R.: Berning, R. F.; and Coffin, L. F., Jr.: High -Temperature, Low-
Cycle Fatigue Behavior of Tantalum. NASA CR-1930, 1971.
21. Kass, J. N.; and Stoner, D. R. : Joining Refractory /Austenitic Bimetal Tubing.
Rep. WANL-PR-(ZZ)-002, Westinghouse Electric Corp. (NASA CR-72353), 1967.
22. Stoner, D. R. : Joining Refractory /Austenitic Bimetal Tubing. Rep. WANL-PR-
(ZZ)-001, Westinghouse Electric Corp. (NASA CR-72275), Oct. 1966.
23. Buckman, R. W. , Jr.; and Goodspeed, R. C. : Evaluation of Refractory /Austenitic
Bimetal Combinations. NASA CR- 15.16.., 1970.
24. Kass, J. N. ; and Stoner, D. R. : Evaluation of Tantalum/316 Stainless Steel Bi-
metallic Tubing. NASA CR-1575, 1970.
25. Thompson, S. R. ; Marble, J. D. ; and Ekvall, R. A. : Development of Optimum
Fabrication Techniques for Brazed Tantalum/Type 316 Stainless Steel Tubular
Transition Joints. Rep. GESP-521, General Electric Co. (NASA CR-72746),
1970.
26. Friedman, Gerald I. : Co-Extruted Tantalum-316 Stainless Steel Bimetallic Joints
and Tubing. Rep. NM-9904.10, Whittaker Corp. (NASA CR-72761), 1970.
27. Stoner, D. R. : Evaluation of Tantalum/316 Stainless Steel Transition Joints. Rep.
WANL-M-FR-72-006, Westinghouse Electric Corp. (NASA CR-12111), Dec. 1972.
45
28. Thompson, S, R. : SNAP-8 Refractory Boiler Development Program: Mercury
Thermal Shock Testing of 2-1/2-Inch-Diameter Bimetallic Joints for SNAP-8
Applications. Rep. GESP-587, General Electric Co. (NASA CR-72829), 1970.
46
NASA-Langley, 1973 17
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